AUTOMOTIVE COOLING AIRFLOW CORRELATIONS
by CHRISTOPHER M. ROSEBERRY, B.S. in M.E
A THESIS
IN MECHANICAL ENGINEERING
Submitted to the Graduate Faculty of Texas Tech University in Partial Fulfillment of the Requirements for the Degree of MASTER OF SCIENCE
IN MECHANICAL ENGINEERING
Approved
August, 1990 ri u ACKNOWLEDGEMENTS
First of all I would like to thank Dr. Walt 01er,the chairman of my committee.. It was my great fortune to work with such a gifted investigator and skilled advisor. Dr. Duane Jordan, who also has served on my graduate committee, I hold in high regard not only for his substantial knowledge and intellect but also for his great dedication to his profession. The third member of my committee. Dr. Tim Maxwell, whom I admire for his seemingly boundless enthusiasm, has contributed significantly to the success of this research through his excellent work as my instructor in Automotive Aerodynamics. I would like to express my appreciation to Ford Motor Company for the generous support, which has gone well beyond that stricUy monetary, and for the permission to publish the findings of this study. I also hold much gratitude for the vitally important counseling I received from Dr. Delores Mack and Dr. Gary Ferrell during my undergraduate studies. Most of all, I thank my wife and my family for the love and support that has made everything possible for me. TABLE OF CONTENTS
ACKNOWLEDGEMENTS ii TABLE OF CONTENTS iii ABSTRACT iv LIST OF FIGURES v NOMENCLATURE vii CHAPTER 1 BACKGROUND 1 CHAPTER 2 TECHNICAL APPROACH 8 2.1 Theoretical Basis 8 2.1.1 Approximate Aerodynamic Analysis 8 2.1.2 Irreversibility Analysis 10 2.2 Test Set-Up 12 2.3 Test Procedure 29 CHAPTER 3 RESULTS 36 CHAPTER 4 CONCLUSIONS AND RECOMMEDATIONS 57 4.1 Conclusions 57 4.2 Recommendations for Further Study 58
REFERENCES 59 APPENDIX A DATA AQUISrrON PROGRAM 60
APPENDIX B MEASUREMENT SYSTEM EVALUATION PROGRAM 66
m ABSTRACT
The design of more efficient automotive cooling systems requires better understanding of the losses in fluid energy through the front end cooling openings. A wind tunnel study with simplified models has been conducted to investigate the factors effecting the cooling opening energy losses. The data from this study have been successfully correlated using newly defined dimensionless parameters. The influence of the studied geometrical factors is clearly identifiable with the data presented in dimensionless form. An algorithm has been devised which can accurately predict the behavior of a combination of cooling openings given the behavior of each of the openings operating individually.
IV LIST OF FIGURES
Figure 1.1 Large Above-Bumper Opening 2 Figure 1.2 Below-Bumper Opening 3 Figure 1.3 Cooling Air Streamtube for Stationary Vehicle with Fan Running 5 Figure 1.4 Cooling Air Streamtube for Vehicle Moving at High Speed 6
Figure 2.1 Cooling Airflow Streamtube 9 Figure 2.2 Front End Portion of Automobile Represented by Model 13 Figure 2.3 Front End Model Attached to Duct 15 Figure 2.4 Front End Model Detached from Duct 16 Figure 2.5 Blower Connected to Duct 17 Figure 2.6 Laminar Flow Element 18 Figure 2.7 Open Test Section with Experimental Apparatus 19 Figure 2.8 Streamline and Bluff Front End Models 20 Figure 2.9 Bluff Model 21 Figure 2.10 Streamline Model 22 Figure 2.11 Disassembled Front End Showing Honeycomb and Screen 25 Figure 2.12 Hose Connection 26
Figure 2.13 Groundplane Supports 27 Figure 2.14 Blower Assembly 28 Figure 2.15 Kiel and Pitot-Static Probe Locations 30
Figure 2.16 Pressure Acquisition Set-Up 31
Figure 2.17 Validyne System 32
Figure 2.18 Experimental Measurements and Calculations 35
Figure 3.1 Total Grille Loss vs. Flow Rate for Streamline Model 37
ff Figure 3.2 Ram Recovery Coefficient vs. Flow Rate for Streamline Model 38
Figure 3.3 Ram Recovery Coefficient vs. Velocity Ratio for Streamline Model... 39 Figure 3.4 Dimensional Data from Schaub and Charles Study 40 Figure 3.5 Dimensionless Data from Schaub and Charles Study 41 Figure 3.6 Frontal Opening Data 42 Figure 3.7 Correlated Frontal Opening Data 44 Figure 3.8 Streamline Model Frontal Opening Data with Correlation 45 Figure 3.9 Bluff Frontal Opening Data with Correlation 46
Figure 3.10 Effect of Varying Bottom Opening Length 48 Figure 3.11 Correlated Bottom Data for Various Bottom Opening Widths 49 Figure 3.12 Effect of Varying Airdam Height 50 Figure 3.13 Effect of Varying Groundplane Clearance on Bottom Opening 51 Figure 3.14 Effect of Varying Groundplane Clearance on Frontal Opening 53 Figure 3.15 Combination Data with Predictions 55 Figure 3.16 Range of Velocity Ratio for Taurus and Cougar 56
VI NOMENCLATURE
P. Freestream Static Pressure u. Freestream Velocity
AP „ Grille Pressure Loss grille
'^loM Grille Loss Coefficient kgriuc Ram Recovery Coefficient
In- Irreversibility Rate wk. Power other than Row Power ex Exergy rii Mass Flow Rate
T. Freestream Temperature To Environmental Temperature S Constant Pressure Specific He k Ratio of Specific Heats R Gas Constant P Density Ma Mach Number p Experimentally Measured Pres: measured V Experimentally Measured Volti meafured v^. Reference Voltage vu V^^ Reference Voltage for Zero Differential Pressure P^ Reference Pressure U model Average Velocity inside Model Uft-ee Freestream Velocity Uop«, Average Velocity through Opening Qmodei Volumetric Flow Rate through Model ^modd Cross-sectional Area of Model ^op« ^^^ of Opening Vlll CHAPTER 1 BACKGROUND Efficiency and reliability have become increasingly important issues in automotive design. The problem of optimally incorporating the cooling system into the overall design often brings these two concerns into conflict. Adequate cooling is essential to provide reliable operation. However, the sizable power requirements to provide convective heat transfer (typically 5 to 10% of total engine output) need to be minimized in the interests of fuel economy. The energy input to the cooling system typically comes from two immediate sources, the radiator fan and ram airflow due to the forward motion of the vehicle. Accordingly, to optimize the cooling system both the external and internal flow conditions must be considered. Moreover, the external and internal flow fields must interact to some degree. The interfaces between the two flow fields are the external cooling openings of the vehicle and the underside of the engine bay. The energy change, due to a loss in total pressure, across the front end cooling openings (which is the focus of this study) is the first in a series of energy changes through the cooling airflow circuit. The other energy changes are due to the effects of the radiator, condenser, radiator fan and shroud, and engine blockage. In addition to being functionally important, the cooling air inlets are very often a distinctive styling feature of a particular make or model automobile. Consequently, there is great variation in the geometry and location of frontal cooling openings. At one extreme are the large vertical grilles located above the bumper which are typical of many of today's luxury cars and the vast majority of past car designs (Figure 1.1). In contrast are the below bumper inlets with litUe or no projected frontal area which are no longer only common to sports cars (Figure 1.2). Since the mid-1970s, the design of more efficient cooling systems has been addressed by several automotive researchers. Olson (1976) recognized the need to introduce scientific methods to replace trial and error methodology in optimizing vehicle cooling systems. By performing full-scale wind tunnel tests with traversing anemometers mounted behind the radiator, Olson identified the effects of various front- end configurations. However, one of the limitations of this study was the necessity of estimating grille airflow by flow visualization. 1 ff F- Cooling Air -^- — Inlet Figure 1.1 Large Above-Bumper Opening »:::j- r- Cooling Air Inlet Figure 1.2 Below-Bumper Opening Hawes (1976) outlined many possible improvements to cooling systems and recognized that an overly large frontal intake will not necessarily increase cooling airflow. Schaub and Charles (1980) direcUy addressed the interaction between the ram airflow and the radiator fan. The streamtube concept was used to present a key argument in the study concerning the front grille loss. The researchers asserted that the entrance area to the cooling airflow streamtube is infinitely large when the vehicle is stationary with the fan running (Figure 1.3); whereas, at high speeds this entrance area can be quite small (less than one square foot) (Figure 1.4). Schaub and Charles offered this extreme streamtube variation as an explanation for the large changes in the pressure losses through the front grille. Like Olson, Schaub and Charles used traversing anemometers mounted behind the radiator which in this study measured velocities and flow rates during both road tests and full-scale wind tunnel tests. To investigate the influence of the fan in the wind tunnel tests, the engine bay backpressure was independently controlled by a catchment sealed to the engine bay underside. This catchment was, in turn, connected by a duct to a separate fan. This approach provided a way to measure the front grille pressure loss over a wide range of cooling flow rates at a constant wind tunnel speed. Williams (1985) showed that grille open area, the amount of grille area that can be frontally projected onto the radiator, is not a reliable parameter to be used in predicting cooling drag and ram airflow. By testing in both aerodynamic and environmental wind tunnels, Williams also determined that over 95% of the physical variation in measured cooling performance (compared to radiator test stand data) could be attributed to factors involved in the cooling airflow. Renn and Gilhaus (1986) placed emphasis on the importance of providing ducting between the cooling openings and the radiator and condenser. This feature increases the effectiveness of the cooling system by preventing the air from bypassing the heat exchangers and by reducing hot air recirculation. These researchers noted that, based on cooling drag data, not many cooling systems were closely optimized. After qualitatively outlining other front-end effects, they concluded "aerodynamic improvements need not necessarily interfere with cooling requirements." oo Streamtube Boundary Streamtube Inlet Area Figure 1.3 Cooling Air Streamtube for Stationary Vehicle with Fan Running Streamtube JStreamtube Boundary Inlet Area Figure 1.4 Cooling Air Streamtube for Vehicle Moving at High Speed [• r^:rr--:r,:r:::-'- A primary difficulty in optimizing automotive cooling systems is the inability to accurately predict the cooling performance in a new vehicle before a production prototype is constructed and tested. Because many production tooling decisions have already been made at this point in the development, cooling system changes can only'be made at great expense. For this reason, cooling systems are frequendy designed conservatively. To address this problem, a computational cooling system model, TTU_COOL, is under development at Texas Tech University by Drs. D. P. Jordan and J. W. Oler with the support and cooperation of Ford Motor Company. Since the interaction of the ram and radiator fan effects on the grille pressure loss has been poorly understood, the application of TTU_COOL has been restricted by the marginal accuracy to which the grille pressure loss can be estimated (especially in the absence of applicable wind tunnel data). This limitation reduces the utility of TTU_COOL in new vehicle development where it offers the most potential benefit. The purpose of this study is to provide a better understanding of the general energy loss behavior of various configurations of frontal cooling openings over a wide range of flow rates and vehicle speeds. The findings of this study are intended to not only be relevantt o TTU_COOL, but also to any design effort involving frontal cooling openings. CHAPTER 2 TECHNICAL APPROACH 2.1 Theoretical Basis 2.1.1 Approximate Aerodynamic Analysis Consider a streamtube passing through an automobile cooling opening (Figure 2.1). The Bernoulli equation P- + /2pUi = P, + KpUf + AP^, (2.1) with a non-recoverable pressure loss, AP^^, may be applied to this situation. APg^ue shall be defined as the Grille Pressure Loss. (In the usage of this study, the word "grille" will refer to any cooling opening without regard to the actual presence of a grating structure.) The Grille Pressure Loss is simply the difference in total pressure across the cooling opening. AP^c = (P. + /2pU-')-(P, + /2pU,^) (2.2) Dividing the equation through by the freestream dynamic pressure gives AP^_j P.-P--t->^pUf KpUr2i ~ * KpU\/^T[2i . (2.3) The left-hand quantity shall be defined as the Grille Loss Coefficient ,k,^, . _ *-"8riUe KpUl (2.4) The equation (2.3) may be rearranged to yield P,..^+)^puf -1 '^l.—lot. y^p^- . (2.5) This arrangement brings the quantity 1 - kj^ together which shall be defined as the Ram Recovery Coefficient, k^^^, k =l-k KgriUe A ^los» (2.6) 8 Figure 2.1 Cooling Airflow Streamtube 2.1.2 Irreversibility Analysis A highly informative means of evaluating the various sources of inefficiency in a particular system is to determine the rate of irreversibility for each process in the system. The irreversibility rate for the grille airflow can be shown to be related to ^Pgriue* kio« ' ^d kg^^ by the following derivation. Consider the inlet of the grille airflow streamtube to be at state «« and the outlet to be at state 1. The conservation of mass principle requires lii^ = riii = rh. The irreversibility rate may be expressed as In- = wk,-Hrii(ex.-i->^Ui)-m(ex,-i-KUf)- -^^ V dt Aystem (2.7) where wkj is power other than flow power from the surroundings, ex is exergy, and 0 is the extensive availability. By observing that only flow power is present in the streamtube and assuming steady- state conditions Eq(2.7) may be simplified to — = (ex.-h/2Ui)-(ex, + KUf) m (2.9) The exergy may be defined as ex = (h - TQS) - (h - Ts)o (2 JQ) where the subscript 0 indicates the properties are evaluated at the most stable state for the substance in equilibrium with the environment. Substituting for ex in Eq(2.9) gives — = [(h - ToS)« - (h - Ts)o + /zUi] - [(h - ToS), - (h - Ts)o + KUf] m . (2.11) This equation may be combined and rearranged to give ^ = (h.-h,)-To(s^-s,)-HK(Ui-Uf) m (2.12) Assuming that air may be treated as an ideal gas with constant specific heats, the changes in enthalpy and entropy may be expressed as h.-h, = c,(T.-T,) ^2.13) 10 and s« - s, = -c . T» k P„ (2.14) ForT,-T.«T. andP,-P^«P^ T» T^ (2.15) and P P -P P_ P_ (2.16) Using these approximations and substituting for h^ - hj and s^ - s, in Eq(2.12) gives T^-T^ k-lfi-P, 2 TT2> —-Cp(T^-T,)-hToC, + /2(U1-U0 m '^ ' T« k P^ (2.17) Noting that TQ = T^ , Eq(2.17) reducest o = -T.c/^]^^ + M(Ui - Uf) 111 (2.18) kR Since (2.19) then m- (2.20) Equation (2.17) may now be expressed as ^ = -^(P,-PJ + /2(Ui-Uf) m P_ (2.21) Note that by the ideal gas equation P, P- = T.R (2.22) Since Ma « 0.3, the flow may be assumed to be incompressible, so that P- = Pi = P. (2.23) Accordingly, Eq(2.21) becomes 11 — = -(P.-p,)+K(ui-uf) m P . (2.24) From Eq(2.24) it may be deduced that Irr _ AP^e m (2.25) Thus, the irreversibility per unit mass is approximately equal to the loss of flow energy per unit mass to the environment Recalling that -(P--Pi)+/2(Ui-u,1 .^ ,^v T2 TT2o> >^U-. , (2.26) it can be concluded that — «k,«.(/2Ul) m , (2.26) or m "^ . (2.27) The irreversibility is clearly dependent on both the internal and external flow. Note that this analysis does not require the flow to be isothermal. 2.2 Test Set-Up The primary objective of this study was to determine the relationships between the pressure losses across front-end cooling openings and the cooling air flow rate, the vehicle speed, and the size and location of the openings. To more clearly reveal these relationships, the design of the experimental wind tunnel models incorporated only the front end portion of an automobile (Figure 2.2). This approach was taken to isolate the front end cooling openings from any unique aerodynamic effects associated with the other geometrical features of a realautomobile . The wind tunnel models consisted of two generically shaped automotive front ends of approximately one-fifth scale. These simplified front-end models were fitted 12 Represented by Model Not Represented by Model Figure 2.2 Front End Portion of Automobile Represented by Model 13 individually to a smooth elongated duct of the same cross-section (Figures 2.3 and 2.4). The duct was, in turn, connected by a hose to the suction side of a high-pressure centrifugal blower (Figure 2.5). The volumetric flowrate through the model was measured by a laminar flow element (Meriam model 50MC2) at the blower exhaust (Figure 2.6). The groundplane was represented by a flat plate mounted beneath the model, which produced a boundary layer much thinner than that of the floor of the wind tunnel. This experiment utilized the open test section of the Texas Tech wind tunnel (Figure 2.7). The ouUet of the nozzle of the wind tunnel is 3 ft. high and 4 ft. wide. The maximum air velocity at the nozzle exit is approximately 78 ft./s. Adjustable vanes at the inlet to the wind tunnel blower provide a means to vary the wind tunnel velocity. Tests were conducted with two geometrically different front end models (Figure 2.8). Much like the front end bodywork of a real car, the inside contours of both models followed the external contours to form a shell. The first model, which will be referred to as the bluff front end, was simply a rectangular prism with slighdy rounded edges (Figure 2.9). In contrast, the second model, which will be called the streamline front end, was a streamlined shape formed from elliptical sections (Figure 2.10). This variation in streamlining reflects the extremes observed in automotive design. To compare the performance of different cooling opening configurations, both models featured frontal openings, bottom openings, and airdams that could be interchanged or eliminated. For both front ends, three frontal opening sizes were used. The openings projected 5%, 10%, and 20% of the model frontal area onto a vertical plane normal to the freestream direction. Although the projected areas were identical, the absolute areas of the frontal openings were slightly larger for the streamline front end than for the corresponding openings on the bluff front end. Each frontal opening had the same height to width aspect ratio as the front-end cross section (1:1.85) and was centered on the longitudinal axis of the model. The bottom openings, which had no projected frontal area, were varied in both length and width (Table 2.1) for a total of five different sizes. In all cases the front edge of the bottom opening was 0.75 in. behind the extreme front of the model. Three different airdams were used that varied in height downward from the model (Table 2.1). Whenever an airdam was in place, it was located at the back edge of the bottom opening. 14 Figure 2.3 Front End Model Attached to Duct 15 Figure 2.4 Front End Model Detached from Duct 16 Figure 2.5 Blower Connected to Duct 17 Figure 2.6 Laminar Flow Element 18 Figure 2.7 Open Test Section with Experimental Apparatus 19 Figure 2.8 Streamline and Bluff Front End Models 20 5" Front + View 13 1/2" r T" 1" 6" Bottom View 1' -X_ 0.25" H i 0.75" 0.5" T — —Longitudinal Axis Side Rounded Edges are 0.25" in Radius View _L^ 0.75" Figure 2.9 Bluff Model 21 '—-I Front View + 5" r—! Bottom View Intersecting Edge Rounded to 1/4" Radius Longitudinal Axis Side Rounded Edges are View 1/4" in Radius Fi gure 2.10 Streamline Model 22 -,-^~r,^,T ,fyf-ir-. ,,-.^.T^T,^.^/»,,,»/»y»^»;Y-f*Y*v*if*J*, Table 2.1 Varied Geometrical Parameters SmaU Nominal Large Frontal 5.06% of 9.95% of * 20.39% of Openings Frontal Area Frontal Area Frontal Area Frontal Area ht.=1.031 in. ht.=1.406 in. ht.=2.031 in. = 0.4 sq. ft. wt.=2.828 in. wt.=4.078 in. wt.=5.781 in. Short Nominal Long Bottom Opening lt.=0.375 in. lt.=0.641 in. lt.=0.968 in. Length wt.=11.50 in. wt.=11.50 in. wt.=11.50 in. Narrow Nominal Wide Bottom Opening lt.=0.641 in. lt.=0.641 in. lt=0.641 in. Width wt.=3.50 in. wt.=7.50 in. wt.=11.50 in. Small Nominal Large Airdam ht.=0.375 in. ht.=0.641 in. ht. =0.969 in. Height wt.=11.50 in. wt.=11.50 in. wt.=11.50 in. Low Nominal High Groundplane gpc=0.925 in. gpc=1.850 in. gpc=2.775 in. Clearance 23 The duct, to which the front end model was affixed, was made of plywood panels (0.25 in. thick) attached to wooden comer members (0.5 in. by 0.75 in.) and was supported by four streamline steel legs. To monitor the average static pressure inside the duct, a ring of small holes was drilled around its interior. The holes were connected by a slot on each side of the duct. Each slot was capped by a thin aluminum plate glued to the outside of the duct. Small brass tubes provided pressure taps to each of the four slots. To obtain steady pressure measurements inside the duct, honeycomb (0.25 in. wide cells, 2 in. long) and finely meshed screen (36 mesh, 59% open area) were placed directly behind the front end model (Figure 2.11). The screen and honeycomb eliminated the large-scale three-dimensional turbulent eddies in the internal flow by dampening both the axial and transverse turbulent components. The static pressure ports were located 8 in. downstream of the honeycomb and screen to allow complete dampening of the turbulence to occur. To connect the hose, the rear of the duct was fitted with a rectangular to round transition (Figure 2.12). However, the hose and the rear of the duct were likely sources of internal and external flow disturbances. To prevent the disturbances from interfering with measurements, the duct was extended beyond the distance to which the perturbations could propagate upstream. The groundplane was constructed of 0.75 in. thick plywood with the leading edge cut to a 45-degree edge. The groundplane restedo n steel angles that clamped to the legs of the model. Additional support to the front of the groundplane was provided by legs clamped to the sides of the test section (Figure 2.13). These supports allowed the groundplane to be adjusted up and down relative to the model. The groundplane clearance or the distance between the bottom of the model and the groundplane, was another geometrical parameter that was varied in the study. In general, three different groundplane clearances were used (Table 2.1). The groundplane extended far in front and to each side of the front end to minimize any unrealistic effects or interference. The groundplane was carefully leveled after every change in groundplane clearance. The entire blower assembly was mounted on a cart (Figure 2.14). The blower impeller was belt-driven by an electric motor (Baldor 1.0 Hp DC), with a speed controller (U.S. Controls). A shon length of rubber hose connected the blower housing with the laminar flow element. This compact arrangement was made feasible 24 Figure 2.11 Disassembled Fron. End Showing Honeycomb and Screen 25 Figure 2.12 Hose Connection 26 Figure 2.13 Groundplane Supports 27 Figure 2.14 Blower Assembly 28 by the inherent insensitivity of the laminar flow element to both upstream and downstream flow disturbances. The laminar flow element features an approximately linear relationship between pressure drop and flow rate. The entire airflow circuit was carefully sealed to prevent any flow measurement errors resulting from infiltration or leakage. Three pressure measurements were needed to determine each data point. The static pressures from each of the four banks of pressure ports on each side of the duct were manifolded together to obtain an average static pressure inside the duct. The freestream total pressure was sensed with a kiel probe located above and in front of the model (Figure 2.15), and measured differentially with respect to the static pressure inside the duct A pitot-static probe, mounted to one side of the kiel probe, was used to sense the freestream dynamic pressure. The pressure drop across the laminar flow element was simply measured using the taps provided on the instrument All pressure data were acquired with a Validyne pressure measurement system connected to a personal computer (UCR with Intel 8086 and 8087 chips) via an Analog to Digital converter (Metrabyte DASH 16) (Figure 2.16). The Validyne system consisted of differential pressure transducers (DP45) (Figure 2.17) connected to a signal processor which output a DC voltage proportional to the differential pressure. The DC voltages, which represented the magnitude and polarity of the applied pressure, were input to the A to D converter. 2.3 Test Procedure The wind tunnel tests were conducted interactively with a data acquisition program (Appendix A). Before the program was initiated, the experimental apparatus was operated at the maximum flowratean d wind tunnel speed with the Validyne system monitoring the pressures. Because the maximum A to D converter range was -5 to +5 volts, the output of each of the three Validyne channels was adjusted to approximately 4.5 volts. This procedure prevented over range voltages while maintaining good measurement resolution. After initiating the data acquisition program, the first step was to recordth e date and time of the test in the data file. The dimensions of the front opening, bottom opening, and the airdam were also recorded followed by the groundplane clearance. 29 Figure 2.15 Kiel and Pitot-Static Probe Locations 30 -. ^ji.^^xiwmrrfmrmmmmmtUtMt Figure 2.16 Pressure Acquisition Set-Up 31 Figure 2.17 Validyne System 32 A zero differential pressure was then applied to the three Validyne transducers and the output from each transducer was adjusted (with a potentiometer on the signal conditioner) to approximately zero volts and recorded into the data file. The program determined all voltage values by averaging 1000 sampled integer values. A reference pressure, generated across the laminar flow element, was simultaneously applied to each of the pressure transducers and an inclined manometer. The manometer reading was typed into the computer which automatically triggered the recording of the corresponding voltage from the transducer. The output voltages from the Validyne system are linearly related to the pressures applied to the transducers. Thus, with a known voltage corresponding to a reference pressure and a voltage for zero differential pressure, any subsequent voltage (within the range of the A to D converter) can be relatedt o a measured pressure from V -V P mcMured zero (P«f.) meafured ref. zoo (2.27) After making the pressure connections with the transducers, the wind tunnel and blower were activated to begin data acquisition. The environmental pressure was read from a barometer and entered into the program. The ambient temperature was entered for the first data point which triggered the recording of the measured pressures. At this point, the program processed the measured data to obtain the freestream velocity (U^), the average air velocity inside the model (U„^^), and the total pressure loss across the front end model (AP^^). With the air density known from the environmental temperature and pressure (making the ideal gas assumption), the freestream velocity was determined by a simple algebraic manipulation. The internal velocity, which will be referred to as the model velocity, was determined by dividing the volumetric flowrate by the internal cross- sectional area (0.4 sq. ft.). The flowratewa s determined from the pressure drop across the laminar flow element using the linear calibration curve provided for the instrument. The calibration curve was incorporated into the program for this purpose. Temperature corrections to the flow rate were made using a curve fit of tabulated temperature correction factors. 33 To determine the front end total pressure loss, the dynamic pressure of the internal flow and the pressure drop due to the honeycomb and screen must be calculated and combined with the measured pressure loss (Figure 2.18). A curve fit of experimentally determined pressure drop versus flowrate data for the particular honeycomb and screen combination was used by the program to calculate this pressure drop. The pressure drop due to the flow straighteners was ordinarily less than the measured pressure drop through the model by approximately an order of magnitude. The above data processing routine was repeatedfo r all subsequent data points in a test run. The quantities recorded into the data file for each data point were the following: total grille pressure drop, the freestream velocity, the model velocity, the freestream dynamic pressure, the internal dynamic pressure, the air density, the temperature, and the measured drop in pressure through the model. The intermediary data were recordeds o that the pertinent parameters (AP^rine,, U^ and U„„^) could be recovered if an error existed in the algorithm or if a systematic error occurred during a test. While the data were acquired automatically (i.e., after each temperature input), the operating point of the experimental apparatus was changed manually. The operating point could be altered by adjusting either the wind tunnel speed or the blower speed. The tests were conducted more effectively by making numerous blower speed adjustments at three to four set wind tunnel speeds. After each data point was determined, pertinent information concerning the data point was displayed on the screen to facilitate the management of a test. Periodically, between experimental test runs, the experimental pressure measurement system (consisting of the computer, A to D converter, and Validyne instruments) was checked with an evaluation program (Appendix B). By comparing the readings from an inclined manometer with the measurements obtained utilizing the experimental system, the evaluation program was able to assess the accuracy of the system. The techniques used for the system checks were consistent with those used for the experimental test runs. 34 p« u« Honeycomb Screen model u, u model W W To^ L Blower Kiel Probe r y^ A (P. + MpUi)-P^ Measured AP.grill^e = (P- + KpUl) - (P, + HpUf) Pj — Pfflodel "*" ^^honeycomb.jcreen u, = u„«,^ APgrille = (P- + >^2pUi) - (P^odel + AP,,„^,,„b.scrce„ + MpULel) \————V/ Calculated Figure 2.18 Experimental Measurements and Calculations 35 CHAPTER 3 RESULTS In the analysis of the experimental data, the use of dimensionless parameters offers great advantage in formulating correlations and identifying trends. Figure 3.1 is a dimensional plot of the loss in total pressure across the grille openings versus flow rate for the streamlined model with nominal frontal opening, bottom opening, airdam, and groundplane clearance. These data are typical in that each freestream velocity results in a separate curve. The objective for forming plots of dimensionless variables is that the number of independent parameters is reduced and data such as that in Figure 3.1 may be represented as a single curve. If the grille pressure loss is expressed as the dimensionless Ram Recovery Coefficient, as in Eq(2.6), Figure 3.2 is obtained. The curves now have a common intercept with the vertical axis, but are still separate. If, in addition, the flow rate is made dimensionless by using the ratio of the average velocity through the model cross section to the freestream velocity, (U„„„,.i/U.), the results presented in Figure 3.3 are obtained. As can be seen, all data collapse onto a single curve thereby indicating a successful correlation of total pressure loss, flow rate and freestream velocity. This collapse occurred for all test configurations when plotted with the above dimensionless parameters. Since the use of dimensionless parameters collapses the data onto a single curve, different symbols will be used in the following figures to designate different configurations instead of ftieestream velocities. As noted in Chapter 1, Schaub and Charles studied the front end pressure loss. Figure 3.4, a dimensional plot of front end pressure loss versus flow rate, is an excerpt from their paper. If the data taken from this figure are presented in dimensionless form. Figure 3.5 is obtained. The use of dimensionless parameters U,^d/U„ and ^grille seems to effectively collapse this full-scale data. This result provides some assurance that observations for the simplified front end models are also applicable to full-scale automobiles. Figure 3.6 shows the pressure drop versus flow rate results, in dimensionless form, for both front ends for all sizes of frontal openings without bottom openings. With decreasing opening area, the pressure losses clearly increase more rapidly with velocity ratio or flow rate. The curves corresponding to the streamline model generally 36 Total Grill Pressurg Loss 101— n 4-) ^r n « n D o •^ + if) 6- n n D 4 DB' -f 4- -I- + + + 2^ --H+ I 0 0 2 4 Q (cubic ft/s) FrGestr9c.-n Velocity (ft/s) 50< + <60< o <70< • <80 StrGGmiinG - Nominal Frontal end Bottom Openings Frontal GoGning Arga/Frontal ATGC = iO/» Figure 3.1 Total Grille Loss vs. Row Rate for Streamline Model 37 Ram RGCovgry CoefficiGnt 0.5^ ""^-^^i ^D D •^ + ^ D 0 XL O D -0.5 + -1 0 4 Q (cubic ft/s) FrGGStrGcm Vglocity (ft/s) 50< + <60< o <70< • <80 StrGOiTilinQ - Nominal Frontal and Bottom OpGnings Frontal Opening ArGo/Frontal Area = 10/1 Figure 3.2 Ram Recovery Coefficient vs.Row Rate for Streamline Model 38 Ram Recovery Coefficient ir 0.5-a +, ^^otpn ^^q>: + on + n o. u 0 uO CD -0.5 1 1 -1 0 0.02 0.04 0.06 0.08 0.1 Umodel/Ufree FreGstream Velocity (ft/s) 50< + <60< o <70< n <80 StrGamlinG - Nominal Frontal and Bottom Openings Frontal Opening Area/Frontal Area = lOX Figure 3.3 Ram Recovery Coefficient vs. Velocity Ratio for Streamline Model 39 (.1 o _J o z g COCL/NG AJR nXNV RATE:0.(mVs) Figure 3.4 Dimensional Data ft-om Schaub and Charles Study 40 Schaub and Charles Data ir- n n n n Gh -f^ sx. • r -1 ^ X -1 n CD X -2 X -3 L __j J 0.05 0.1 0.15 0.2 UmodQl/Ufree Vghicle spged (km/h) 1976 Granada X 51 o 103 Figure 3.5 Dimensionless Data from Schaub and Charles Study 41 Ram RecovQry CoefficiGnt +J + 0.5 w^ %^ vvo^ f «i % 0 CD ct 5^ EI s 1 -0.5 V i!f X -r 1_^ i V 0 0.02 0.04 0.06 0.08 0.1 Umodel/Ufree Frontal Openings - Percent of Frontal Area Bluff a 5% o 10A + 20>: Streamline x 5X v 10% A 202 Figure 3.6 Frontal Opening Data 42 follow those of the bluff front end but depart at low velocity ratios to lower intercepts with the vertical axis. From the geometry of the bluff front end, it is reasonable to expect that at zero velocity ratio or flow rate a frontal opening will produce the freestream stagnation pressure behind the opening which corresponds to a Ram Recovery Coefficient of one. For the streamline front end, the average pressure coefficient over the opening area, at zero velocity ratio, would be expected to be less than unity because the opening is curved and somewhat oblique to the freestream direction. The corresponding Ram Recovery Coefficients at zero velocity ratio are expected to be less than that of the bluff model and less than unity. The selection of the area used in calculating a velocity from the flow rate through the model is an arbitrary consideration. If the area of the opening is used to define the velocity (U^ = Q^^/A^) and U^ is used instead of U^^^in the velocity ratio. Figure 3.7 is obtained from the same data presented in Figure 3.6. As can be observed, using U„p„ in the dimensionless velocity successfully correlates the data for the various frontal openings. A correlation is easily obtained for the dimensionless data by noting the slope and vertical intercept of a plot of ln(ki^^) versus ln(U<^/U_). k^ = l-2.4596(U^/U.f (3 J) Figure 3.8 shows the correlation with the streamline data while Figure 3.9 shows the bluff data with the same correlation. In Figure 3.8 as in Figure 3.6, the streamline model produces Ram Recovery Coefficients of less than unity for zero velocity ratio. In Figures 3.7, 3.8, and 3.9, only data in the range of 0.01 < U^JU. < 0.1 have been plotted so that all data are in the same range of flow rates. However, it should be noted that the same U^,p„/U^ for different opening areas does not correspond to the same flow rate (for a given freestream velocity). In regards to the bottom opening, the bluff and the streamline models yield distinctly different results. Flow visualization conducted with the Bluff model revealed a large separation bubble which blanketed the bottom opening. The Reynolds number, which was approximately 500,000 based on model width, was not sufficient to keep the flow attached past a 0.25 in. radius edge. As a result of the opening being located in a region of low pressure separated flow, the Ram Recovery Coefficients generally 43 Ram Recovery Coefficient ^'-^:. I 0.5 L -1 CD 0 0.2 0. 4 0. 6 0.8 Uopen/Ufree Frontal Openings - Percent of Frontal Area Bluff ° 5;^ ^ 10/^ + 2o;{ Streamline x 5X v 107. A 20X Figure 3.7 Correlated Frontal Opening Data 44 „,.,.,.,.,-,.,••:' Ram Recovery Coefficient CD -0.5- 0. 4 0. 6 Uopen/Ufree Streamline Front End - Percent of Frontal Area D 5% + lo;: A 20/; kGrill = 1 - 2.4596*(Uopen/Ufree) Figure 3.8 Streamline Model Frontal Opening Data with Correlation 45 til:-'':';'''- ' • . -.-.-vT .J • J1JI.I.L1 .L'l^wwii in., mn I •• Ram Recovery Coefficient c CD 0. 4 0. 6 0.8 Uopen/Ufree Bluff Front End - Percent of Frontal Area • 5% + 10/. A 2QZ kGrill = 1 - 2.4596*(Uopen/Ufree) Figure 3.9 Bluff Frontal Opening Data with Correlation 46 began at less than zero and progressively decreased with velocity ratio. Under this condition it was common for the flow inside the model to reverse direction and flow out of the opening with the blower running at low speed. This situation is not characteristic of actual vehicles and consequendy the data for the bottom openings on the bluff model are not presented. Figure 3.10 illustrates the effect of varying the bottom opening length on the streamline model. The opening was lengthened by moving the airdam and rear edge of the opening back from the front edge of the opening. The widths of the opening and airdam were kept constant. Using U„p^ in the dimensionless velocity ratio does not collapse the curves as it did with the frontal openings. However, the curves do follow the same trend and show a linear decrease in Ram Recovery Coefficient with velocity ratio. Surprisingly, the Ram Recovery Coefficient decreases with increasing opening area. This trend is believed to be a result of increasing the distance of the airdam from the front edge of the model. As the airdam is moved back, the average pressure over the opening decreases for a given velocity ratio, thereby reducing the Ram Recovery Coefficient The effect of varying the bottom opening width is shown in Figure 3.11. The airdam dimensions were held constant as was the length of the opening. The data for the different widths are reasonably correlated by k^ = 0.37 - L827(U^/U J ^3 ^^ The fixed airdam position apparently causes the Ram Recovery Coefficient to be approximately the same for each opening width at a particular velocity ratio. Figure 3.12 shows that ram recovery is improved by increasing the airdam height. The dimensions of the opening were held constant while the height of the airdam was increasingly extended down from the model. Because the opening area has been held constant, there is no reason for using U„p„ instead of U^^^ in this figure. As with the other data for bottom openings, the Ram Recovery Coefficient exhibits a linear dependence on velocity ratio. Decreasing the groundplane clearance improves ram recovery for a bottom opening (Figure 3.13). The size and position of the opening and airdam were fixed 47 Ram Recovery Coefficient T~^ CD 0.4 0.6 Uopen/Ufree Bottom Openings - Opening Length/Model Width Streamline x .0278 v .0475 A .0718 Figure 3.10 Effect of Varying Bottom Opening Length 48 Ram Recovery Coefficient T T 1 7- n 0.5 CD -O.Sr- 0.6 Uopen/Ufree Bottom Openings - Opening Width/Model Width Streamline n . 259 o .555 + .852 - kGri11=0.37-1. 827(Uopen/Ufree) Figure 3.11 Correlated Bottom Data for Various Bottom Opening Widths 49 Ram Recoverv Coefficient CD 0.04 0.06 0.1 Umodel/Ufree Airdam Height/Groundplane Clearence Streamline • .203 o .346 + .524 Figure 3.12 Effect of Varying Airdam Height 50 Rom Recovery CoefficiQnt 0 ^^^ ^%fa^$ O m R3 D •++ ^ O O CD + % o + OO D -1 -H- O — + + D + -2 1 0 0.02 0.04 0.06 0.08 0.1 Umodel/Ufree Groundplane Clearence/Model Height Streamline • . 184 o .368 + .552 Figure 3.13 Effect of Varying Groundplane Clearence on Bottom Opening 51 while the groundplane clearance was varied. The degree of influence of the groundplane clearly increases as it approaches the model. The curve fits in Figure 3.14 reveal that groundplane clearance has a minimal effect on frontal opening Ram Recovery Coefficients for both the bluff and the streamline models. For the streamline model, the groundplane clearance seems to influence the proportion of the airflow going under the model to that over the model. Extensive data were collected for configurations that combined frontal and bottom openings. The same geometrical parameters were varied for the combination of openings as for the individual openings. The combination data generally yields linear curves with a reduced slope relative to the data for the isolated bottom openings. The factors that produce higher Ram Recovery Coefficients for the bottom opening and the frontal opening have the same effect for the combinations of openings. An algorithm was devised to predict the Ram Recovery Coefficients for a particular combination of openings using the data from the individual openings tested separately. The primary assumption for this method is that the drop in total pressure is the same for both the frontal and bottom opening flow paths. Accordingly, the Grille Loss Coefficients for the two openings must match since the upstream conditions are identical for both openings. Treating the Ram Recovery Coefficient as the independent variable, the corresponding velocity ratios (U„p„/U..) are found from correlations for each opening. 1/B U open X uope n = -(C-k ghiii U. D ic) Jbooom (3.5) Both velocity ratios are multiplied by Uie ratio of the opening area to the cross-sectional area and added together to obtain the velocity ratio (U„,^/U^) for the combination for the particular Ram Recovery Coefficient U model uope n open uope n open A u« u. Jcambiiution u. - fronUlL model J Jbooom ^ model (3.6) 52 Ram Recovery Coefficient r— CD 0.02 0. 04 0. 06 0.1 Umodel/Ufree Groundplane Clearence/Front Height Bluff --B-0.184 -^0.368 -4-0.552 Streamline _^o^ 194 ^,5^0.368 -^0.552 Figure 3.14 Effect of Varying Groundplane Clearence on Frontal Opening 53 Figure 3.15 shows data from combined configurations and predictions from the data for separately tested openings. The accuracy of the predictions validate the assumption made for the algorithm. To demonstrate that the data collected in this study are in a range relevant to real automobiles, data from full-scale wind tunnel tests are presented in Figure 3.16. U„K>dei/U« has been plotted against flow rate for a 1989 Cougar and 1987 Taurus. An area of 10 ft.^ was used to obtain U„,^ from the flow rate. Based on this figure, the range of interest for the velocity ratio is observed to be 0.03 to 0.1. The data from this study, which ranges from 0.01 to 0.2, completely covers the range of interest. 54 Ram Recovery Coefficient "^^n r L. 0 0.02 0.04 0.06 0.08 Umodel/UfrGQ StreamlinG - Front Openings With Nominal Bottom Opening Predictions from Separate Data Sets Data from Combined Configuration x 5Z • lOZ + 20Z Figure 3.15 Combination Data with Predictions 55 ^ ^^ Umodel/Ufree vs. Coolinq Flow Rate a Of 0.1^ • + Of <> V + O 0.05- ^v + A n + A^ n ° >^^ n 1 i i i 1 1 i } 0^- _l-. - ..., _..J . 1 „_ J 0 1000 2000 3000 4000 Q (cfm) Vehicle speed (mphmph) 1989 Cougar • 30 o 45 + 75 1987 Taurus X 30 V 60 A 90 Figure 3.16 Range of Velocity Ratio for Taurus and Cougar 56 CHAPTER 4 CONCLUSIONS AND RECOMMENDATIONS 4.1 Conclusions The dimensionless parameters defined in this study have greatly facilitated the understanding of the behavior of front end cooling openings by reducing the number of required variables from three to two for any front end configuration. The Ram Recovery Coefficient and a velocity ratio based on an arbitrarily defined internal velocity and the freestream velocity replace the dimensional parameters of grille pressure loss, freestream velocity, and cooling flow rate. The performance of the isolated frontal openings is correlated by the relation k^. = l-2.4596(U^/UJ^ (4.1) where U is the average velocity through the cooling opening. Ram recovery for a bottom opening may be improved by increasing the height of the airdam, decreasing the groundplane clearance, and decreasing the distance of the airdam from the leading edge of a vehicle. The correlation for a bottom opening with the airdam at a nominal distance from the leading edge is k^ = 0.37-1.827(U^/U.) ^^ 2) The performance of a combination of cooling openings can be accurately predicted with correlations from the individual openings of that combination. The primary assumption is that the Ram Recovery Coefficient must be the same for each opening of a combination at a particular operating point. Using the correlations for frontal openings and bottom openings, a velocity ratio can be determined for the combination of openings by the relation U model uope n open uope n open A + u. A u. combination u. JfionulL model J J bottom .^modcl J (4.3) 57 4.2 Recommendations for Further Study The influence of various grille gratings on the Ram Recovery Coefficient needs to be investigated in experiments with a Reynolds number at or close to that of full scale. Because the openings in this study were either normal, almost normal, or parallel to the freestream direction, the behavior of oblique openings warrants further study. A better understanding of bottom openings may be obtained by varying the length of the opening while the airdam is fixed in position relative to the leading edge of the model or vehicle. Varying the area of the bottom opening while maintaining a constant aspect ratio may also prove to be enlightening. If the technique used is sufficienUy non-invasive, determining the pressure coefficient distribution of the external and internal surfaces of the front end may reveal the flow mechanisms that control the Ram Recovery Coefficient. 58 REFERENCES Hawes, S.P. "Improved Passenger Car Cooling Systems." SAE Paper 760112. February, 1976. Olson, M.E. "Aerodynamic Effects of Front End Design on Automobile Engine Cooling Systems." SAE Paper 760188. February, 1976. Renn, V. and A. Gilhaus. "Aerodynamics of Vehicle Cooling Systems." Journal of Wind Engineering and Industrial Aerodynamics. Vol 22:339-346. 1986. Schaub, U.W. and H.N. Charles. "Ram Air Effects on the Air Side Cooling System Performance of a Typical North American Passenger Car." SAE Paper 800032. February, 1980. Williams, J. "An Automotive Front-End Design Approach for Improved Aerodynamics and Cooling." SAE Paper 850281. 1985. 59 APPENDIX A DATA AQUISmON PROGRAM r frend.c Chris Rosebenry July 89 data aquisition program for thesis experiment V #include "stdio.h" #include "io.h" #include "dos.h" #include "math.h" #include "grphx.h" #include "fcntl.h" #include "stat.h" #include "dacq.h" void config(void); void set_zeroes(void); void set_spans(void): void get_data(void); void Delay(void); float lntoVolt(long int.int); FILE *output: struct expar { float pModel; float qFree; float pLFE: }; long int sum; int 1; struct expar zero.Pspan.Vspan.data; void main() ( configO: set_zeroes(); set_spans(); get_dataO: } void configO { r input date, time, model configuration V struct date ( int month; int day; int year; }: struct time { int hour; int minutes; ); 60 stmctdate d; stmcttlme t; float foh,fow.lol,low,adamht,gpc; output=fopen("exp.dat","w"); printf("\n\n\n Enter date (mm/dd/yr)-"); scanf("%d/%d/%d".&d.month.&d.day.&d.year); fprintf(output."%d/%d/%d",d.month,d.day,d.year); printf("\n\n Enter time of day (hh:mm)-"); scanf("%d:%d",&t.hour,&t.minutes); fprintf(output."\n%d:%d",t.hour.t.minutes); printf("\n\n Enter front opening height (Inches) ="); scanf("%r.&foh); fprintf(output,"\n %f".foh); printf("\n\n Enter front opening width (inches) ="); scanf("%r,&fow); fprintf(output,"\n %f",fow); printf("\n\n Enter lower opening length (Inches) ="); scanf("%f".&loO; fprintf(output,"\n%fMol); printf("\n\n Enter lower opening width (inches) ="); scanf("%f",&low); fprintf(output,"\n %f",tow); printf("\n\n Enter airdam height (inches) ="); scanf("%r,&adamht); fprintf(output,"\n %f",adamht): printf("\n\n Enter groundplane clearence (inches) ="); scanf("%f",&gpc); fprintf (output,"\n %f",gpc); g_clear(); } void set_zeroes() { r get voltages for zero pressure gradients on transducers */ d_open(16,0x310.2,4); initCounter(10); printf("\n\n\n Set zeroes on Validyne"); getcharO; printf("\n\n Hit 61 sum=a_d1(3)+sum; DelayO; } zero.qFree=lntoVolt(sum,i); fprintf (output,"\n %f",zero.qFree); printf("\n Zero for qFree = %f volts",zero.qFree); getchart); printf("\n\n Hit 62 printf("%f volts for qFree".Vspan.qFree); printf("\n\n Enter span pressure for pLFE (in. H20) ="); scanf (-%f".&Pspan.pLFE); getchart); printf("\n\n Hit V while(T>0.0) { sum1=0; sum2=0; sum3=0; for(i=1;l<=1000;+-H) { sum1=a_d1(2)+sum1; sum2=a_d1 (3)+sum2; sum3=a_d1 (4)+sum3; DelayO; } data.pModel=lntoVolt(sum1 ,i); data.qFree=lntoVolt(sum2.i); data.pLFE=lntoVolt(sum3,i); Tcorr=(-0.8034E-7)*rrT+(0.2686E-4)*rT-(0.5903E-2)*T+1.309; r Pcon'= Patm/29.92; V pModel=((data.pModel-zero.pModel)/(Vspan.pModel-zero.pModel)) •Pspan.pModePconv; T b / sq ft V 63 MpModel=pModel/conv; /* In. H20 V pLFE=((data.pLFE-zero.pLFE)/(Vspan.pLFE-zero.pLFE))*Pspan.pLFE; rin.H20V Qcfm=(405.0/7.7)'pLFE; Qacfm=Qcfm*Tcon^ Umodel= QacfnVO.4/60.0; r ft/sec */ Patmc=Patm* 144/2.036; rho = (Patmc/(53.34*(459.67 + T)))/32.174; qModel=Umodel*Umoderrho/2.0; pHS=-(0.0000001606)*Umoder Umoder Umodel+(0.0009413)*Umoder Umodel +(0.003424)*Umodel-(0.0001757);/* in. H20 V pHS=(pHS/27.71)*144;/* b/sqftV qFree=((data.qFree-zero.qFree)/(Vspan.qFree-zero.qFree))*Pspan.qFree*conv; MqFree = qFree/conv; T in. H20 V Ufree=sqrt((qFree*2.0/rho)); pGriU= pModel - pHS -qModel; If (qFree > 0.0001) NpGrill=pGrill/qFree; else NpGrilU 10000; if (Ufree> 0.0001) Uratio= Umodel/Ufree; else Uratlo = 10000; g_clearO; printf("\n\n\n Total pressure loss = %f (Ib/sq ft)",pGrilI); printf("\n\n\n Wind tunnel speed = %f (ft/s)",Ufree); printf("\n\n\n Air velocity in model = %f (ft/s)",Umodel); printf("\n\n\n Uratio = %f, pGrill/qFree = %f ",Uratio,NpGrill); printf("\n pModel = %f in. H20. qFree = %f in. H20, pLFE = %f in.H20",MpModel,MqFree,pLFE); printf("\n\n\n Want to record datapoint? (1-y,2-n)"); scanf(-%d",&ichs); if (ichs != 2) { fprintf(output,"\n %f %f %f %f %f %f %f %r,pGrill,Ufree,Umodel, qFree,qModel,rho,TpModel); } getcharO; printf("\n\n Enter temperature (F) for next datapoint="); scanf ("%r.&T); getcharO; } d_close(); pGrilU 1010.0; fprintf(output,"\n %f %f %f %f %f %f %f",pGrill,Ufree,Umodel, qFree,qModel,rtK).T); fclose(output); } void DelayO ( r initCounter(100);V r waitQ; V } 64 float lntoVolt(long int Sum ,int I) { float dat.datV; dat=((Sum/l)/1.0)-2047.5; datV=(dat/2047.5)*5.0; retum(datV); } 65 APPENDDC B MEASUREMENT SYSTEM EVALUATION PROGRAM /* test.c program for checking accuracy of Validyne transducers */ #include "dacq.h" #include "stdio.h" #include "io.h" #include "dos.h" #include "math.h" #include "fcntl.h" #include "stat.h" float lntoVolt(long int.int); tong int sum; float zero.spanvolt.spanpress,press1,press2.voltpress2.perdis,perful; int chan.i; void main() { printf("\n\n Program to test accuracy of Validyne transducers"); printf("\n\n Input Dash-16 channel to which the transducer is connected "); scanf ("%d",&chan); if ((Chan > 7) || (chan < 0)) exit(1); d_open(16,0x310,2.7); initCounter(IOO); printf("\n\n Adjust Validyne for zero reading"); getcharO; printf("\n\n Hit 66 t>eginning of while loop V whiie(press1 < 9.5) { printf("\n\n Apply, then key in artjitrary pressure within set range (10 to quit)"); printf("\n Pressurel ="); scanf("%f",&press1); getcharO; printf("\n Hit 67 PERMISSION TO COPY In presenting this thesis in partial fulfillment of the requirements for a master's degree at Texas Tech University, I agree that the Library and my major department shall make it freely avail able for research purposes. Permission to copy this thesis for scholarly purposes may be granted by the Director of the Library or my major professor. It is understood that any copying or publication of this thesis for financial gain shall not be allowed without my further written permission and that any user may be liable for copy right infringement. Disagree (Permission not granted) Agree (Permission granted) Jin/&>L, Student's signature Student>s signature ^ -7/3/ /9V Date Date