**Notes on Mark VI Web Model**

This is the first official release of the Mark VI Web version of the
DOE/ORNL HPDM, which is a major upgrade to the Mark V Web model. The new
Web model uses the Mark VI calculation engine, Version 6.1, attached to a
revised Web interface that provides access to improved HX and fan modeling
capabilities.

Many of the improvements are discussed in a new
Powerpoint presentation downloadable from the Mark VI home page. In the
latest Web version, we have implemented new HX-related capabilities of the Mark
VI engine such as improved airflow modeling options, condenser circuit merging,
additional air-side HX surfaces, flow configuration effects, compressor voltage
effects, and new parametrics. The following describes the new capabilities of
the Mark VI Web model
relative to the existing official Mark V version.

**Improved HFC Thermodynamic and Transport Properties,
Flow Control and Heat Transfer Correlations**

New thermodynamic properties have been incorporated for HFC R410A and R-404A for superheated, two-phase, and subcooled refrigerant regions. Previously, in the Mark V (PUREZ) model, the subcooled properties were obtained from saturated properties at the same temperature. The new online presentation shows the improvement in COP predictions from the improved R-410A property representations. The improvements in R-410A and R-404A properties are most significant at condensing temperatures between 130F and the critical temperatures just above 160F.

We also recently updated the transport properties (viscosity, thermal conductivity, and specific heat calculations) for R-410A, R-404A, R-507A based on work of Geller (2000). The most significant change was in the liquid viscosity values for R-410A which are much lower now than those used in Mark V. We also revised the liquid viscosity values for R-22 downward based on new REFPROP 6 values (1999). Plots of these corrections to the viscosity values for R-410A and R-22 are shown in the new online presentation as well.

New flow control correlations for R-410A in short-tube orifices and capillary tubes were added to the Mark VI engine. The R-410A-specific short-tube orifice correlation is from Payne and O'Neal (1999). The R-410A capillary tube correlation is an adaptation by Rice (2001) of an R-410A-specific correlation by Wolf et al, 1995 to correct for improved R-410A viscosity properties. In the Mark V model, an R-12/R-22 based capillary tube correlation from ASHRAE (1975 Equipment Volume) is used for performance predictions with HFCs. We have not adopted the generalized capillary tube correlation by Bittle et al (1998) recommended by ASHRAE (1998 Refrigeration Volume) because it was found to have been developed with the same overestimated liquid viscosity properties for R-410A as in the work of Wolf (1995).

The evaporation heat transfer correlation has been updated to use the more HFC-suitable correlations of Bivens and Yokozeki (1994) with a dryout point specified at 90% quality. This correlation is an adaptation of those developed by Wattelet et al (1993) and Jung (1989). The condensation heat transfer correlation was also updated to use a more HFC-suitable correlation of Dobson (1994) for annular and wavy-stratified flow. The average two-phase heat transfer coefficients for both evaporation and condensation are now determined for a heat flux assumed to vary with local overall air-to-refrigerant conductance. The programming structure of the algorithm calculating the average evaporator and condenser heat transfer coefficients was also modified to make it easier to update to newer correlations in the future. The new correlations give about 25% lower condensing coefficients and 35% lower evaporating coefficients than with Mark V for R-410A.

We added a new option to specify rifled tubes which will increase the
effective refrigerant-side heat transfer coefficient by 90% and the pressure
drop by 50%. One can use the user-specified refrigerant-side multipliers to
adjust these values, if needed.

**New Airflow Calculation Options**

In the Mark V and earlier models, the specified volumetric airflow rates for the indoor and outdoor coils were the actual rates (acfms) based on the inlet conditions.In Mark VI, we have changed this to improve the way the air side mass flow rates are calculated from given volumetric flows. New ways of specifying and adjusting airflow as inlet conditions change have been provided.

Both new approaches start by having the airflow specified at standard conditions (scfm). This is essentially specifying a given air mass flow rate.

To generate manufacturers' performance tables, this air mass flow is usually fixed for a range of indoor dry and wet bulb conditions. The model now maintains this fixed scfm if the fan reference temperature option is not selected (by leaving the fan reference temperature box unchecked).

The second option is to check the fan reference temp. box and provide the fan reference temperature at which the given volumetric airflow (scfm) was obtained. If the box is checked but no value is given, default values are provided suitable for draw through fans in cooling or heating mode. With this information, the model will calculate the fan characteristic airflow capability and will adjust the mass flow rate of air through the fan inversely proportional to the fan inlet temperature.

In this way, this second approach models the change in mass flow as the inlet
conditions to the fan change. This is especially significant for heat pumps,
where the decrease in indoor mass flow from cooling to heating mode can be as
much as 10%. The same options are provided for the outdoor fan where the mass
flow increases between cooling to heating mode (hotter to cooler air over the
fan) can be more than 20%.

**Indoor Air Handler Options to Specify Filter and
Heater Sizes**

Users can now specify the filter area directly or by giving the applied nominal capacity of the filter. The default filter size is for a capacity of 2.5 tons cooling so if a 5-ton capacity is specified, the filter area will be automatically doubled. (Previously a fixed filter size was assumed.)

For heat pump operation, the number of heater banks and the cross-sectional
area of the heater section can now also be specified.These settings will
affect performance predictions only if the fan power is determined from first
principles using calculated pressure drops. However, they will also affect the
indoor blower/fan efficiency levels deduced for a given fan power as given in
the detailed text output under the fan/blower performance section.

**New Air-Side Fin Patterns and Heat Transfer
Correlations**

The number of available air-side fin patterns has been increased from three (smooth, corrugated, and slit) to six (smooth, corrugated, and slit, louvered, convex louvered, and smooth wavy). The air-side heat transfer correlations for all the surfaces are new and are based on the work of C.C. Wang and associates at the Industrial Technology Research Institute (ITRI) in Taiwan. (See reference list below.) Each correlation is an empirical function of coil geometry and airflow. Examples of the predicted heat transfer coefficients as a function of face velocity for indoor and outdoor coils (as evaporators and condensers) are shown in the new online presentation, where the predictions of the Mark V model are shown as dotted lines. (The Mark V correlations were based on a smooth fin correlations of Gray and Webb (1986) with simple multipliers of 1.45 and 1.75 used for corrugated and slit fin surfaces, respectively.)

Fin surface type has also been added as a parametric so that all the fin surfaces can be compared using default fin pattern geometries or patterns specified by the user. For the five patterned fin surfaces, default fin patterns have been set in the model. However, the user can modify these by using the fin pattern check boxes on the indoor and outdoor coil pages. By selecting the pattern graphic, a more detailed pattern geometry graphic is displayed, showing the input geometric parameters for the selected surface. (The fin patterns graphics were adapted from Fig.1 of Wang, C.C., 2000. "Recent Progress on the Air-Side Performance of Fin-and-Tube Heat Exchangers", International Journal of Heat Exchangers, 1:49-76.

If smooth fins are selected and the fin pattern box is checked on the indoor and outdoor coil pages, any of the default pattern settings are available to be changed. This is provided for use when fin type parametrics are to be run, allowing fin patterns to be set for separately for each surface, if desired. (Otherwise, default fin geometry settings will be used.)

New correlations have also been included to correct dry heat transfer coefficients for wet surface conditions. These are also shown in the Mark VI online presentation along with the previous smooth fin correction used in Mark V based on the work of Myers.

Corresponding new air-side pressure drop correlations have not yet been added to the model. As such, tradeoffs of air-side heat transfer vs pressure drop effects should not be made yet with this model until this has been completed. We hope this will be accomplished in the next few months. The model is currently using the air-side pressure drop correlations from Mark V for smooth, wavy (corrugated or smooth wavy), or enhanced (slit, louvered, and convex louvered) surfaces.

For the first time, the ORNL model now calculates separately the added
air-side surface for certain patterned fins (for corrugated, convex louvered,
and smooth wavy fins). The calculated air-side multipliers can be found in the
detailed text output under the "Computed Heat Exchanger Characteristics" output.
The user can adjust these calculated air-side area multipliers if needed in the
indoor and outdoor unit multipliers section (any user provided multipliers are
applied to the internally calculated ones). We have assumed that there is no
area multiplier for louvered or slit/lanced surfaces, but they can be provided
here as well, if desired.

**Condenser Circuit Merging and Flow Configuration
Capability**

Users can now specify circuit merging for the outdoor or the indoor coil when used as a condenser. This allows the condenser subcooled region to have fewer parallel circuits than in the two-phase region, which is often the case in outdoor coils designed for AC-only operation.

The HX arrangement can now be given as N-row crossflow, cross-counterflow, or cross-parallelflow. In the Mark V model, only the crossflow option is provided. This option controls how the heat transfer in the single-phase region is calculated using different effectiveness()-NTU relationships for the different 1- to N-row configurations. The proper -NTU relationship for refrigerant-side mixed, air-side unmixed is selected depending on the specified configuration and the number of rows. (For the two-phase regions with pure, azeotropic, or near-azeotropic refrigerants, the -NTU equation is flow configuration independent.)

Historically, the DOE/ORNL HPDM has always assumed a crossflow arrangement
with each refrigerant region seeing the same inlet air temperature, as described
in the Mark V
online presentation. However, in the Mark VI model, for the
cross-counterflow arrangement in the condenser, we assume that the subcooling
region is ahead of the two-phase region and that the superheated region is
behind the two-phase. The number of rows occupied by each single-phase region is
calculated and the appropriate 1-N row equation is applied. (We have not yet
made a similar configuration change for evaporators or for the
cross-parallelflow configuration. Such changes are planned as part of
modifications planned next year to model Hx performance with gliding
refrigerants. ) For flow configurations other than cross-counterflow condensers,
all refrigerant regions are presently assumed to see the air at the same
entering temperature as is the case throughout in the Mark V Web model. In the
Mark V model, the single-phase region of each HX is also modeled with a many-row
crossflow -NTU relationship assuming both refrigerant and air streams are
unmixed.

**New Line Heat Loss Options**

On the Refrigerant Lines page, the heat losses in the discharge and liquid
lines can now alternatively be given as temperature changes, just as has been
the case for the suction line.

**Additional Liquid Volumes**

Means were added to specify additional volumes where liquid refrigerant can
reside. These are for the liquid/filter dryer and for the outdoor and indoor
liquid headers. These are after the Refrigerant Lines section under Component
Dimensions, the same place where the accumulator and compressor free volumes are
specified. By providing good estimates for these volumes, the absolute and trend
accuracy of the charge inventory balance model should be improved.

**New Parametric Ambient Control
Options**

One can now specify the compressor performance calibration factors for power
and/or mass flow rate as a function of ambient in cooling or heating mode. Line
heat losses or gains (or temperature changes) can also be given as a function of
ambient. (Later, we hope to add an option for calculating line heat losses or
gains within the program.) In the heating mode, one can also add a heating load
line versus ambient so that the model can calculate the required backup heat and
overall COP including resistance heat as a function of ambient.

**New Compressor Motor Performance Estimates and
Accounting for Compressor Voltage Effects**

In the Mark VI model, we have upgraded the model of compressor motor
efficiency and speed as functions of torque and voltage for a baseline 230V
single-speed 1-phase motor. The new efficiency and speed curves as a function of
nominal operating torque and for different voltages are shown in the new online
Powerpoint presentation. This provides more accurate estimates of the motor
operating efficiency and speed than with the three-phase motor model in Mark V
and allows a compressor map to be modified to account for over- or under-voltage
operation from 230V. This capability is of value in correcting for voltage
effects in lab tests when the nominal voltage was not maintained and also in the
evaluation of the maximum power draw of a compressor under peak conditions at
reduced voltage.

**Further Model Plans**

We plan to add a few more parametrics to the Mark VI Web version soon. This includes compressor EER and capacity scaling factors and requires that we make a transition to scaling compressor EER and capacity scaling within the core model rather than externally in the Web script at present.

Our near term plans are to add a compressor database with compressor map
coefficients from one or more compressor manufacturers. Also we intend to add
the new air-side pressure drop correlations consistent with the recently
upgraded air-side heat transfer correlations. Work is planned in the coming year
to complete changes needed to accommodate gliding refrigerants such as R-407C.
These latter two efforts are a part of work begun in Sept. 2002 by Purdue and
ORNL related to ASHRAE Techincal
Research Project 1173 (TRP-1173) on improvement and validation of AC design
models for use with HFC refrigerants.

**Status of Mark V Web Model and Data
Sets**

Please note that we intend to maintain the Mark V Web model at its current site indefinitely for users who do not have a pressing need to change over to the Mark VI version at the present time. As the heat exchanger air- and refrigerant-side correlations have changed significantly from the Mark V version, new HX calibrations are recommended when switching from Mark V validated data sets to Mark VI.

Web data sets created and saved with the Mark V model will need to be edited
to run with the Mark VI model. This can be done by making the changes described
at
this link. In the next few months, we plan to provide a more convenient way
to convert these files by a batch process.

**REFERENCES**

**Refrigerant-Side Heat Transfer Correlations**

**Evaporating**

Bivens, D. B., Yokozeki, A., "Heat Transfer Coefficients and Transport Properties for Alternative Refrigerants", International Refrigeration Conference Proceedings, Purdue Univ., 1994.

**Condensing**

Dobson, M. K. et al., Heat Transfer and Flow Regimes During Condensation in
Horizontal Tubes, ACRC TR-57, Univ. of Illinois, May 1994, 242 pp.

**Refrigerant Transport Properties**

McLinden, M. O., S. A. Klein, E. A. Lemmon, and A. P. Peskin, 1999. *NIST
Thermodynamic and Transport Properties of Refrigerants and Refrigerant
Mixtures--REFPROP, Version 6.01*.

Geller, V. Z., D. Bivens, and A. Yokozeki, 2000. "Viscosity of Mixed
Refrigerants R404A, R407C, R410A, and R507A", *8 ^{th} International
Refrigeration Conference at Purdue University*, West Lafayette, Indiana, USA,
July 25-28, 2000, pp. 399-406.

Geller, V. Z., B. V. Nemzer, and U. V. Cheremnykh, 2000. "Thermal
Conductivity of Mixed Refrigerants", *14 ^{th} Symposium on
Thermophysical Properties*, June 25-30, 2000, Boulder, Colorado.

**Refrigerant Flow Controls**

Wolf, D. A., R. R. Bittle, and M. B. Pate, 1995. *Adiabatic Capillary Tube
Performance with Alternative Refrigerants, ASHRAE RP-762, Final Report*,
Engineering Research Institute, Iowa State University, ERI-95413, May.

Bittle, R. R., D. A. Wolf, and M. B. Pate, 1998. "A Generalized Performance
Prediction Method for Adiabatic Capillary Tubes", *HVAC&R Research*,
Vol. 4, No. 1, January, pp.27-44.

ASHRAE, 1998. *1998 ASHRAE Handbook, Refrigeration*, *IP-Edition*,
pp. 45.24-45.27.

ASHRAE, 1975. *1975 ASHRAE Handbook, Equipment*, Chapter 20, pp.
20.20-20.27.

Payne, W. Vance and D. L. O'Neal, June 1999. "Multiphase Flow of Refrigerant
410A Through Short-Tube Orifices", *ASHRAE Transactions*, Vol. 105, Part 2.

**Air-Side Fin-and-Tube HX Heat Transfer Correlations**

__Smooth Fins__

Wang, C. C., et al, 2000. Heat Transfer And Friction Characteristics Of Plain
Fin-and-Tube Heat Exchangers, Part I: New Experimental Data, *International
Journal Of Heat And Mass Transfer*, Volume: 43, Issue: 15 August 1, 2000, pp.
2681-2691.

Wang, C. C., et al, 2000. Heat Transfer And Friction Characteristics Of Plain
Fin-and-Tube Heat Exchangers, Part II: Correlation, *International Journal Of
Heat And Mass Transfer*, Volume: 43, Issue: 15 August 1, 2000, pp. 2693-2700

__Corrugated Fins__

For tube diameters >= 1/2",

Wang, C. C., Y. T. Lin, C. J. Lee, and Y.
J. Chang, "Investigation of Wavy Fin-and-Tube Heat Exchangers: A Contribution to
Databank", *Experimental Heat Transfer*, 12:73-89(1999)

For tube diameters < 1/2",

Wang, C. C., J. Y. Jang, and N. F. Chiou,
"A Heat Transfer and Friction Correlation For Wavy Fin-and-tube Heat
Exchangers", *International Journal of Heat and Mass Transfer, *Vol
42(1999) pp.1919-1924.

__Slit Fins__

Wang, C. C., 2001, "A Comparative Study of Compact Enhanced Fin-and-Tube Heat
Exchangers", *Int. J. Heat and Mass Transfer*, Vol. 44, pp. 3565-3573.

__Louvered Fins__

Wang, C. C., C. J. Lee, C. T. Chang, and S. P. Lin, "Heat Transfer and
Friction Correlation for Compact Louvered Fin-and-tube Heat Exchangers",
*International Journal of Heat and Mass Transfer*, Vol 42 (1999)
pp.1945-1956.

__Convex Louvered Fins__

Wang, C. C., Y. M. Tsai, and D. C. Lu, "Comprehensive Study of Convex-louver
and Wavy Fin-and-tube Heat Exchangers", *Journal of Thermophysics and Heat
Transfer, *Vol 12, No. 3, July-September 1998, pp.423-430.

__Smooth Wavy Fins__

Mirth, D. R. and S. Ramadhyani, 1994,"Correlations for Predicting the Airside
Nusselt Numbers and Friction Factors in Chilled-water Cooling Coils",
*Experimental Heat Transfer*, Vol 7: 143-162.

__Wet Coil Adjustment
Factor__

Wang, C. C., et al, 2001. "Empirical Airside Correlations of Fin-and-Tube
Heat Exchangers Under Dehumidifying Conditions", *Int. J. of Heat
Exchangers*, Vol. 2, pp. 151-178.

__Summary Air-Side Heat Transfer
Papers__

Wang, C. C., "On the Airside Performance of Fin-and-Tube Heat Exchangers",
*Heat Transfer Enhancement of Heat Exchangers*, S. Kakac, et al. (eds.),
Kluwer Academic Publishers, pp. 141-162, 1999.

Wang, C. C., 2000. Recent Progress on the Air-Side Performance of
Fin-and-Tube Heat Exchangers, *International Journal of Heat Exchangers*,
1:49-76.